under cut and fill
TRANSCRIPT
Introduction
Backfilling in North America has been practiced since the turn of the century. De Souza et al. (2003)
have summarized the advancements in backfill with the introduction of hydraulic fills in the 1950s
and the addition of cement in the 1960s. This, coupled with cemented rock and paste fills being
introduced in the 1980s and 1990s, respectively, resulted in the implementation of mining methods
that require extraction under a consolidated back largely comprised of fill rather than timbered
mats/cables (Marcinyshyn, 1996). The increased use of consolidated fills in the late 1990s to
present, under engineered conditions with a high degree of reproducibility in terms of strengths and
predicted behaviours, has enabled man-entry methods such as underhand cut-and-fill to be
implemented under greater controlled spans,resulting in a safe and economic alternative to
conventional cut-and-fill mining (Mah et al., 2003). A database of twelve underhand cut-and-fill
operations was compiled 12 part of this project through mine visits. The operations were located
throughout North America and summarized in Table 1.
The placement of consolidated fill, either cemented rock fill or paste, requires one to understand the
overall factors affecting design. Figure 1 graphically summarizes some of the parameters that are
being investigated in terms of their implication on developing a design span enabling man entry
access. A sill for this study is defined as a consolidated layer of previously placed fill immediately
above the mine opening that is being excavated. This sill may be comprized of one large vertical
height placed by bulk mining, as shown in Figure 2a, or by single lifts as placed by conventional drift
and fill and/or underhand cutand- fill as shown in Figure 2b. The major difference is that in Figure
2a, one is largely operating remotely from the immediate filled back, whereas in Figure 2b, one is
mining by man-entry methods. This necessitates that the factor of safety for the man entry be
substantially greater than the non-entry approach.
Fig. 1. Mining underneath consolidated backfill
DESIGN CONSTRAINTS
Figure 1 shows the factors that have to be accounted for in terms of mining under an engineered
back. These will be outlined in this paper from a general perspective with a focus on the analytical,
numerical assessment of span and applied loading conditions.
DESIGN LOAD
Fig. 2. Mining under a consolidated back
A critical factor is estimating the design loads onto the sill mat. In a recently completed M.A.Sc.
Thesis, Caceres (2005) employed Placer Dome’s Musselwhite mine as a case study and looked at
the loading conditions that exist on a cemented rock fill mat, as shown schematically in Figure 2a.
Knowing the loads is critical to determining the strength required of the sill mat for the given stope
geometry. Underestimating can cause a premature failure of the sill mat once mining exposes the
mat, whereas overestimating can result in unnecessary expense due to the cost of the cement in
place. Knowing the vertical loading is not a trivial solution, many factors affect the overall loading
conditions, as evidenced from the many theoretical derivations that are available [see Janssen
(1895), Terzaghi et al. (1996), Reimbert (1976) and Blight (1984)], all of which have significant
assumptions in terms of the coefficient of lateral earth pressure “K” as detailed by Marcinyshyn
(1996). The value for “K” describes the ratio between the horizontal and vertical stresses in the fill
and, indirectly the ability of the degree of load transfer by arching. When fill is placed initially, very
little shear resistance is mobilized through grain interaction, the coefficient of earth pressure at rest
(Ko). Subsequent placement of fill causes the fill mass to settle and compact, increasing the shear
resistance and transfer load to the abutments through arching. As mining underneath the sill pillar
progresses, a void is created towards which the fill mass will tend to move if unconsolidated with the
transfer of vertical stresses laterally through arching. This condition is described by the passive
earth pressure coefficient (Kp) where full shear resistance is mobilized. This is analogous to the
classical embankment theory where the walls are moving into the fill. The effect of ‘K’ employing
individual fill load formulae is shown in Figure 3a. The analytical methods shown in Figure 3a
assume vertical stope walls, which is generally a conservative estimate of vertical loading (Caceres,
2005). The typical geometry was modelled employing FLAC2D (Itasca, 2005) which did not have
the constraints that the analytical methods had in terms of ‘K’ and stope inclination. An analytical
approximation, as shown by equation 1, was derived by Caceres (2005) relating the numerical
simulation to an equivalent relationship, as shown in Figure 3b.
The above was derived for cemented rock fill; however, the analytical solution would be similar to
that of paste as the input parameters would define the loading conditions.
Fig. 3. Estimation of vertical loading onto sill mat
FAILURE MECHANISM
The methodology of span design under consolidated fill is complex because many factors control
the overall stability, as shown in Figure 1. The failure modes and combination thereof must be
analyzed with respect to the placed fill, stope geometry, loading conditions, seismic affects, stope
closure, and support placement, as well as other factors that are due to filling practices, such as
cold joints and gaps between successive lifts among others. This paper employs analytical,
numerical, and empirical tools to attempt to provide an initial tool for the operator for design. The
database of underhand stopes observed by the author is shown in Table 1, which is comprised of
12 operations that include seven cemented rock fill and five having paste within the immediate
back.
The unconfined compressive strength is typically the parameter employed to benchmark the overall
stability of the immediate back. Figure 4 shows the compiled database (Table 1) of backfill
unconfined compressive strengths adapted from De Souza et al. (2003). The design methods
(Table 1) employed a form of limit equilibrium analysis coupled with modelling. The failure modes,
summarized by Mitchell (1991), are shown in Figure 5.
Fig. 4. Backfill strength vs binder content as depicted in Table 1.
Flexural instability was found to be most critical in the absence of rotational instability and closure
stresses (sc) which have to be evaluated separately. Stone (1993) had concluded that for cemented
rock fills crushing, caving, and sliding are generally negated when the sill thickness exceeds 0.5
times the span; absence of closure stresses and the unconfined compressive strength of the
cemented rock fill is greater than 1.5 MPa; and rotational instability, where kinematically possible,
has to be analyzed separately. Figure 6 shows the database that has been compiled in Table 1,
plotted onto a stability chart adapted from Stone (1993), and developed for the design of sills with
vertical sidewalls with a factor of safety of two. The chart is based on flexural instability employing
fixed beam analysis with surcharge loading, using equation 1. It shows the unconfined compressive
strength required (FS=2) for a given sill thickness and span exposed and related to actual field
observations.
Fig. 5. Limit equilibrium criteria adapted from Mitchell (1991).
Fig. 6. Stability chart for the design of undercut sills with vertical sidewalls with a FS of 2. Chart is
based on a fixed beam bending failure with surcharge, adapted from Stone (1993).
Generally, the mine data was found to be more conservative than that required for a factor of safety
of 2.0. This may reflect the quality control requirements at individual operations, along with other
factors, such as seismicity and stope geometry, as shown in Figure 1.
The limit equilibrium approach shown in Figure 5 was simulated by Caceres (2005) employing
FLAC2D models (finite difference code) for a given value of cohesion, span, and stope dip. The
cemented rock fill properties assigned are for a Mohr-Coulomb type of material with strain-softening
behaviour where integrity is lost after 1.5% strain (Swan and Brummer, 2001). The resultant mode
of failure was analyzed for 75o to 90o dip stopes with cohesion on the hangingwall contact varying
from zero to maximum (cemented rock fill cohesion), as shown in Figure 7 for the 75o stope dip.
The analytical approach summarized by Mitchell assumes no hangingwall cohesion for the
rotational instability and this was found to result in a high degree of conservatism. The FLAC2D
simulation shows the failure mode that results for a given stope span, sill height, and cohesive
strength for various assumptions on wall friction. The design curves have been developed for
cemented rock fill operations; however, they are largely dependent on material strength
characteristics that are input into the model. This method does not have the constraints associated
with the limit equilibrium techniques shown in Figure 5 because the relationships have been derived
through numerical simulation of modelled conditions. The critical parameters identified in Figure 7
do not incorporate a factor of safety because they are identified in the numerical model by being
either stable or unstable. Therefore, one must apply a safety factor on the input parameters to
ensure that adequate design factor of safety is incorporated. The cemented rock fill database
shown in Table 1 was overlain onto Figure 7. The type of failure occurring is indicated for each
curve by either “S” for sliding, “Rc” for rotational-crushing, “Rb” for rotational-breaking, and “F” for
flexural failure modes. The degree of cohesion at the wall contact allows flexibility in design,
depending on the quality of fill placement. It must be recognized that, based on the cases identified
in Table 1 and plotted onto Figure 7, the depth of mat largely corresponds to between 50% and
100% cohesion on the hangingwall contact.
Fig. 7. Sill mat stability as derived from numerical modelling; factor of safety > 1.0.
OTHER FACTORS
The above attempts to outline a methodology for span design. It is critical that the method be
calibrated for individual sites, incorporating critical factors such as seismic conditions, installed
support, and methods of fill placement, as these all play a significant role in ensuring a safe
exposed operating span. A major benefit of mining under paste is the mitigation of the hazards
posed by bursting (Blake and Hedley, 2003).
SEISMIC CASE HISTORY — MANAGING ROCKBURSTS AT HECLA’S LUCKY FRIDAY MINE, MULLAN, IDAHO
The following case history has been compiled by Blake and Hedley (2003). Its importance is that
the underhand mining, as practiced at Lucky Friday (mine No. 12 in Table 1), is the first to
incorporate paste to mitigate burst damage, and that the method has been adopted at mines
throughout North America, such as the Red Lake mine in Ontario (Mah et al., 2003) and the
Stillwater mine in Montana (Jordan et al., 2003).
Hecla initiated overhand cutand-fill mining on the Silver Vein at the Lucky Friday mine in the late
1950s. By the mid-1960s, mining had progressed down to the 3050 level (~930 m below surface)
and the mining geometry consisted of long, flat-backed stopes all at the same elevation being
carried up from two or more levels simultaneously. A burst prone sill pillar was formed when mining
from below would approach the overlying mined-out level. As a result of a double rockburst fatality
in 1969, the mining front was changed to a “centre lead stope” geometry. In 1973, the first
computer-controlled seismic monitoring system was installed, and pillar distressing was routinely
carried out when a sill pillar was mined to approximately 12 m (thickness).
This rockburst strategy allowed mining to proceed safely down to below the 4660 level (~1420m
below surface). In 1982, the mining front entered a highly burst prone formation, and serious
rockburst problems were encountered. As a result of rockburst fatalities in 1984 and 1985, Hecla
initiated an experimental underhand cut-and-fill stope along the east abutment of the mine. After
another rockburst fatality in March 1986, Hecla realized that it was not possible to manage their
rockburst problem with overhand cut-and-fill mining. Production mining at the Lucky Friday was
stopped in April 1986, and plans were made to convert the entire mine to mechanized underhand
cut-and-fill mining geometry, which they named Lucky Friday underhand longwall (LFUL). The key
features of this mining method were that pillars would never be formed, and the mining would be
carried out under a stable, engineered, paste type backfill.
Production mining at Lucky Friday resumed in October 1987 incorporating the above changes.
Despite increased rates of rockbursting, as well as larger magnitude bursts (M1 4.1), underhand
cutand- fill mining at the Lucky Friday mine has been carried out without any serious rockburst
injuries or fatalities, and with greatly increased productivity at significantly reduced costs.
Underhand mining has allowed Hecla to very effectively manage their rockburst problem. The
miners have a higher sense of security working below an engineered back. Management has said
that the mine would likely have never reopened after 1986 had it not been for all the benefits of
LFUL mining.
Finally, the paste backfill is only very rarely damaged by the effects of nearby rockbursts. The only
burst-induced fill failure at the mine occurred in 1991 during mining of a remnant pillar where a M1
3.5 burst caused the wall to fail and, in turn, undercut the paste back which collapsed. The peak
particle velocity at the hangingwall/fill mat was approximately 1 m/sec. Despite closure from
ongoing mining, as well as closure and shock loading from the burst, the fill was not rubbilized as
might have been expected.
Fig. 8. Lucky Friday mat
OTHER OBSERVATIONS
The focus of the present research is towards mining under paste. However, a much larger database
of information exists for mining under cemented rock fill, as derived from the Nevada database of
mines operating within a weak rock mass (Brady et al., 2005; Table 1). The limit equilibrium
methods are similar whether one is working with a cemented rock beam or a consolidated paste.
The differences lie in the resultant strengths associated with each, as shown in Figure 4, where the
cemented rock fills exhibit strengths (UCS) generally in excess of two to three times that of the
cemented pastes. This is largely due to cement being able to be more evenly mixed with the larger
aggregate as compared to the paste, which is generally between (60% passing) 20 microns (fine) to
100 microns (coarse) for Mines Nos. 1 and 12 in Table 1. The typical cemented rock fill optimum
aggregate size is 50,000 microns or 5 cm (2 in.) for the database analyazed. The cohesion for paste
fills was estimated as being 0.25 times the unconfined compressive strength based on an internal
angle of friction (f) of 30o and derived from the Mohr-Coulomb relationship where the “Unconfined
compressive strength = 2*Cohesion*(cos f)/(1-sin f).” A value of internal angle of friction of 35o to
40o was employed for the rock fills. The tensile strength was generally derived from the unconfined
compressive strength for consolidated fills and a value of “0.1 ? UCS = Tensile Strength” was
employed (Jaeger and Cook, 1976). A Young’s modulus for paste ranged from 0.6 GPa to 3 MPa
(laboratory) for 10% binder. These values were field calibrated (Williams et al., 2004) to reflect field
data through a combination of earth pressure cells embedded within a paste stope at the Lucky
Friday mine (No. 12) and the closure recorded and related by the relationship Stress = Modulus x
Strain, where the strain was measured by means of closure metres divided by the stope width. This
is shown in Figure 9 with the resultant closure/load history. It is interesting to note that upon
reaching 4 MPa (UCS), the loads showed yielding of the paste. The closure at the Lucky Friday was
in excess of 25 mm. The rock mass of the wall contact was under 50% (RMR76). Mine No. 1
showed closure values of under 10 mm at similar depths, with the wall contact having an RMR76 =
75%. Both sites are considered to be burst prone. A further observation by Tesarik et al. (2005) was
that the earth pressure cells are significantly affected by the paste cure temperature, which can
reach 40oC with stresses measured in excess of 69 kPa due to the temperature difference. The
vertical pressures that arise for a 3 m pour height of paste fill would approach fill pressures of 0.02
MPa/m of pour height (SG=2.0) or 60 kPa of vertical pressure, which is largely equivalent to the
temperature correction on the earth pressure cell. The horizontal pressure would be a fraction of
this. It is important to recognize that the correction factor can be equivalent to the absolute value
measured.
CONCLUSIONS
Mining under consolidated fills is becoming competitive to conventional cut-and-fill mining. This is
because increased spans and productivities are realized through reduced placement of ground
support and, due to working under an engineered back, more control on the mine cycle. This
requires a thorough understanding of the mechanism of support that one is relying on, which is the
consolidated fill immediately above. The fill may be supported in terms of conventional bolts and
screen in order to counter “cold joints” that may develop in the fill, account for variability in fill quality
control, and/or increase the overall factor of safety required due to seismic events in the close
proximity. This requires an understanding of the stabilization effect of the consolidated fill and the
mine environment that it is placed within. Through the gathering of site data, modelling of behaviour
either analytically and/or numerically, and observation and measurement, one will be able to
advance the overall design criteria to provide a safe and cost-effective workplace.
Fig. 9. Instrumented underhand stope at Lucky Friday (Williams et al., 2001)
ACKNOWLEDGMENTS
The authors would like to thank Goldcorp, Barrick, Newmont, Placer Dome, Cameco, Hecla, Coeur
d’Alene Mines, Stillwater, and NIOSH for assisting in the overall project and providing direction.
Particular thanks to NSERC, NIOSH, and individual mining companies for providing funding of the
graduate research that allows industry related research to be conducted.